Text

Forwards Data Re Fuel Performance at Min Subcritical Margin & Hot Channel Thermal Hydraulic Analysis for Steamline Break Accident
	ML19220C332
Person / Time
Site:	Crane
Issue date:	11/18/1977
From:	Herbein J; Metropolitan Edison Co
To:	Varga S; Office of Nuclear Reactor Regulation
References
	TASK-TF, TASK-TMR GQL-1593, NUDOCS 7904300425
	Download: ML19220C332 (77)
	v • d • e

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ATTACEMENT I I. Intreduction Appendix 153 of the mI-2 FSAR presents analyses of worst-case steamline break accidents assu=ing single failures in the non-safety grade =itigating equipment. As a result of a series of =eetings between the NRC Staff and GPUSC cubNting in an appeals =eeting on March 29, 1977, we have co=- zitted to provide ssf ety grade =itigating equipment in accordance with the Staff's letter of March 15, 1977. The Staff agreed at the March 29th =eet-ing that the Appendix 153 analysis represented suitable justification for centinued operation of the plant until the modificationa could be made at the first refueling outage, pri=arily because the specific seismically induced accident scenario was highly unlikely. The following analysis responds to additional Staff concerns regarding the Appendix 153 analysis. This new analysis is very conservative with respect to core perfor=ance, principally because of the s1=plified =cdeling of the feedvater and stean syste=s. This bounding analysis de=enstrates that there is no fuel damage resulting fre= a steamline break, as postulated in Appendix 153. The s1=plifi feedwater/ steam syste= =edeling, as well as other assu=ptions, are so conservative that they actually represent an unmitigated steamline break scenario with respect to core perforrsnee. Thus, there is addi-tional quantitative evidence that operation of MI ' until the first re-fueling outage does not represent an undue risk to the health and saf ety of the public for two reasons: First, the unlikaliness of the postulated seisnically induced unmitigated accident, and seccnd, added assurance of acceptable fuel perfor=ance even if the event should occur. k /l s. I-l

II. Description of ore Performance Evaluation Cases A. General Assu=ptions for Accident Analysis A listing of general analysis assumptions is given in Table 1. The assumed double-ending abrupt guillotine rupture of a =ain steas line provides the most rapid depressurization of the steam generators. No credit was taken for the bachpressure provided to the unaffected steam generator by the large volume of steam pipfng between the unaff ected steam generator and the break. Mair. feedvater was assumed to runout to its maxi =us flow rate with no runback or isolation and no degradation as a result of anticipated hydraulic upsets. Emergency feedvater was assured to be initiated very early (two seconds af ter rupture) in order to bound the additional cooling ef fect of e=ergency feedvater inj ection. The effect of all of the above asaucptions is to maximize the rate of decrease of secondary fluid temperature in the steam generators and to maintain the greatest nucleate boiling h' eat transf er surf ace area. Thus, heat transfer in the steam generators, and consequently depressur-1:ation and overcooling of the pri=ary system, are =ax1=12ed. 1 - p p- / (. ... V I-2

The accident was initiated free rated full power with nominal conditions, except as noted. Full power operation maximizes-the initial inventory on the secondary side of the steam generators. The greatest steam generator inventory based on fouled tube conditions was assumed. Full power operation maximizes the feedwater flow rates and steam flow rates since the feedwater and turbine control valves are initially in the most fully open normal operating positions. Full power operation also maximizes the core return to power conditions, since the delayed neutron precursors are at maxi =um equilibrium concentrations. The accident was assu=ed to occur at the end of first cycle opera-tion. The most conservative (largest negative) fuel and moderator temperature reactivity feedback parameters occur at end of core life, resulting in the greatest potential for return to po-er during cool-down. Control rod worth was based on the minimum shutdown rod worth of 2.1 ok/k at hot shutdown conditions with the most reactive control red stuck out. The possibility that the high pressure injection piping may be filled with fluid at a lower boron concentration than the borated water storage tank was bounded by assuming no boron reactivity effect frem high pressure injection prior to the occurance of the mini =um suberitical margin. However, high pressure injaction was assumed to be actuated as expected during the accident, with no boron reactivity effect. The high pressure injection promotes overcooling by injecting cold water and has a minor effect on maintaining pressure in the pri=ary coolant systen, which delays core flooding tank inj ection. jU JL s I-3

Of f site pcwer was assu=ed to be available throughout the accident in order to naximize =ain f eedwater flow. The accident is analyzed, with and without coastdown of the reactor coolant pe=ps, and results will be presented for both cases. The assu=ptions chosen for the systems analysis of this accident, which are. highlighted above, provide the worst credible conditions for overcooling of the reactor =oderator and a potential return to Other cc binations of assenptions regarding break size, pcuer level, power. the operation of plant equipment, etc., have been evaluated and deter =ined to be less severe. The assunption of blevdown of both steam generators instead of one has been centioned by the Staff and will therefore be discussed in detail. 'w'ich continuous f e:dir.g and blowdown cf one or both steas generators, the primary coolant system pressure and te=perature decrease continueusly. In the evaluatien of the stea= line break accident with single failures presented in the FSAR, Appendix 153, and also in the core perf or:ance evaluation cases presented herein, the loss of suberitical reactivity cargin due to cooldown is terminated following actuation of the core flooding tanks. That is, for the first cycle of operation, it has been demonstrated that the reactor coolant system will depressurice to the actuation pressure of the core flooding tanks ((

1g) bef ore the moderator temperature decreases to the point at which a sign-ificant return to power or return to criticality occurs.

In addition, the rate of core flooding (wi*' cn injection) is more than sufficient to terminate the return to power caused by cooldown. The maxi =us return j i li t_ c I-3 a

to power occurs during the fastest cooldown conditions, pri=arily because the delayed neutron precursors are still at high concentratiens whi t. can ccuse an increase in the fission rate when the subcritical =argin is sufficiently small. When the cooldown rate is less rapid, a =uch less severe increase in the fission rate will occur for a given suberitical =argin. Thur, a s= aller break si:e, lower f eedwater flow rates, or stea= isolation wi.1 create a less severe transient. For secondary syste= accidents in which the core flooding tanks are not actuated, or in which the ficcding rate is s=all, the decrease in suberitical =argin due to cocidown will be :::e than cenpensated by Lorca inj ection fre= high pressure injection. (' dish ccncentration boron inj ectien f rc= the high pressure inj ection would occur within one =inute following actuation.) In this analysis, attention was given Oc realistic =cdeling of phenonena which =ight tend to delay syste= depressurization (and core flood tank acevation) while allowing cooldewn of the =oderator. These effects include early initiation of high pressure inj ection, pressuri=er surge flow, stea= flashing in the coolant hot legs, and stes= flashing in the seni-stagnant reactor vessel upper head region. 'Je conclude that the cases analy=ed represent the worst conditions with regard to core return to power. e9 P, O \\ 'j d v - I-4

e O ~ 3. General Assumptions for Cote Thermal Hydraulic Analysis Core performance during the stea= line break accidents presented herein was evaluated using a conservative, si=plified technique. This =ethod was described in detail at a meeting with the Staf f documented in Reference 1. Details of the input and results of the core thermal-hydraulic calculation will be presented in later sections. Ps W _.1 ($ l-w I-5

G O C. Computer Codes Used A listing of the computer codes used in this analysis and the ap;11 cable topical reports is given in Table 2. The RADAR and PDQ07 codes have been reviewed and approved by the Staff. The TRAP 2 code is currently under review by the Staff. The interaction of cenputer codes for this analysis is described as follows. The PDQ07 code was used to determine the reactivity feedback parameters input to TRAP 2, fitted to core average conditions. These pa.ameters were determined in an iterative manner, with T3AP2 supplying the transient core average conditions which were used to establish the PDQ07 model. The steam line break accident analysis was performed using the TRAP 2 computer code. Transient power level, core pressure, core inlet enthalpy, and core ficw rata calculated by TRAP 2 were corrected for control band errors and then input to the RADAR code. The core thermal hydraulic calculation was perfor=ed with RADAR. Radial peaking conditions input to RADAR were determined by an iterative procedure, using PDQ07 to calculate the radial peaking as a function of =oderator voiding in :Sc stuck control rod assembly, and using RADAR to determine the mini =um DNER and channel voiding as a function of radial peaking. Since the TRAP 2 cceputer code is currently under review by the Staff, a brief discussion supporting its credibility will be provided below. TRAP 2 has been used for secondary system accident analysis on a number of plants employing the B&W 177 Puel Assembly and 205 Fuel Asse=bly NSS designs. The code has been benchnarked against the RELAP-3B cceputer code for a 205 FA stesn line break analysis, with good abreemant / [- J t_ v I-6

of results. The calculational methods used to solve conservation of mass, energy, and or.entum are identical to those in the B&P CRAFT cc=puter code, which has bench::u rked well against experi=entcl results. TRAP 2 and CRAFT have been found to compare vell with each other en a stess line break analysis, for the 177 FA design Oconee plant. The core kinetics solution in TRAP 2 has been compared to that in the B&W CADDS code, with good agree =ent of the power transient during a steam line break. The TRAP 2 calculations performed for the Three Mile Island Unit 2 steam line break analysis have been verified with hand calculations of the key results, such as reactivity feedback, heat transfer rates, and the pressure - temperature relationship in the reactor. These and other comparisens shich have been made give reasonable assurance that the TRAP 2 code and user =odeling vill provide an accurate representation of the consequences of secondary system accidents such as the steam line break. P "? / U .v I-7

D. Key Inputs to the TRAP 2 Code A listing of key inpute to the TRAP 2 computer code is provided in '~ Table 3. Nuclear parmeters are based on end of core life, Cycle 1. The moderator boron concentration of 17 ppm is the minimus value occuring at end of life. The moderator temperature coef ficient of reactivity is closely related to the core boron concentration, and beccmes more negative as the boron concentration decreases. This can be seen in Figure 1. The reactivity worth of injected boron, for example from the core flooding tanks, is less for higher boren con-centrations. However, this effect is more than compensated by the more negative mederator coef ficient which occurs with lower boron concentrations. It is possible that portions of the core flooding lines may contain fluid at a lower boren concentration than the core f ?oeding tankt or the borated water storage tank. This has been bounded by assuming that the fluid isolated frem the core flooding tanks by check salves is at the mini =um cold shutdown boron reactivity, based on no xenon and max 1=um worth centrol red stuck out. The curve of reactivity as a function of core average moderator density, input directly to IRAP2, is shown in Figure 1. The curve corresponding to 17 ppm boron concentration was used in this analysis. The calculation of this curve with PDQ07 was keyed to the transient coolant pressure and te=perature and fuel te=peratue occuring during the accident. Figure 2 shows reactivity insertion as a functica of average fuel camperature. .?e effect of decr;astug moderator te=perature on fuel te=percture feedback was accounted f or in the PDQ07 calculations used to generate this curve. / (. j f. 4 I-8

Figure 3 presents the assumed control rod worth as a function of average moderator te=perature. The cc=bination of this curve with that in Figure 1 is equivalent to a curva of reactivity versus moderator density with control rods in. Figure 4 presents the inverse boron worth of injected boron as a function of average moderator temperature. A constant value of 82.5 ppe/ hK/K was used for these analyses. e w ~ t ,(; tV I-9

O E. KEY INPUTS TO THE RADAR CODE The thermal-hydraulic evaluation of the return-to-power following a main steam line break used the closed channel transient RADAR code (BAW-10069A, Rev. 1). Two sub-channels were =cdeled into the RADAR code. The first sub-enannel represents an average core channel and is used to calculate the core pressure drop as a function of time during the transient. The core pressure drop, as calculated by RADAR, agrees with the steady-state codes in which the entire core 's modeled. The input to RADAR was calculated using the TRAP code and consists of system flow, core power, system pressure, and coclant inlet tc=perature as a function of time. The core pressure drop versus time was applied across a second channel to calculate the minimum DNBR, c;ad temperature, fuel temperature, etc. This second channel assumed the worst Nuclear, Ther=al, and Mechanical conditions exiat s1=ultaneously. If this particular channel meets al.1 the ther=al design criteria, then it can be safely assu=ed that all other channels in the core will be no worse ther= ally than this limiting ch annel. Since the core pressure drop versus ti=e is the only driving force for coolant ficw in the hot channel, any increased heating, voiding, etc. ot.ly decreases the calculated channel flowrate. The =axi=um design conditions are represented by the follcwing assu=ptions: The limiting fuel asse=bly (radial peak of 1.69) as shown in Figure 5 possesses a. the fuel pin with the =ax1=um value of F H (1.78) as deter =ined f rom exa= . tion of the reference design radial peaking distribution and from examination of the =axi=um, nc=inal, and minimum fuel asse=bly spacing to find the effects on the local peaking distribution. b. The =axt=um value of F nuclear (=ax/ avg axial fuel red heat input) is deter =ined for the 11=iting steady-state condition such that sufficient =argin is provided to acce==odate asy==etrical axial power shape considera:io s ,9 / (' ~ 4-I-10

O uhich will permit core operation sith both positive and negative power imbalance. The maximum errors on core pressure and inlet temperature are applied in the most conse rvative manner. They are -65 psia and +2F, respectively. c. Every channel in the core is assumed to have the nominal pressure drop sssociated with the core flow conditions. d. The limiting fuel assembly is penalized 5% in isothermal flow such that it receives only 95% of the flow associated with an average core bundle at 100% of system flow (four-pump operation). The limiting fuel assembly is assumed to have a reduced peripheral flow e. area due to the proximity of the adjacent fuel asse:blies, f. On the subchannel types with maxi =um values of F E aH' channel factors are applied: 1. The channel is assu=ed to have a reduced flow area, represented by w the hot channel f actor F,, less than 1.0. 2. The fuel pin will have the greatest heat output by virtue of the F ", 9 which increases the local surface heat flux. 3. And F, which increases the overall fuel pin power rating, will be greater than 1.0. M ,U .Ju I-ll

O O III. Analysis Results A. Results of the TRAP 2 Analysis Tuc cases were analyzed. In the first case, the reactor coolant pumps were assumed to oparate nornally throughout the transient. Full coolant flev maximi:es heat transf er in the steam generators and thus maxinizes moderator cooldown and the potential for a return to power. In the second case, all four reactor coolant pu=ps were assumed to begin coastdown coincident with rupture of a =ain steam line. Coolant flow coastdown produces a less severe potential for return to pcwer, but is adverse with regards *.o minimum DN3R in the s'utk control rod assembly. The results of both cases will be presented below. 1. Full Reactor Coolant Flow Case A chrononological sequence of events for this case is given in Table 4, along with the sequence of events for the reactor coolant flew ceasedcun case. The response of key system variables is shcun as a function of tine in Figures 6 through 11. Note that rupture occurs at 1 second in Figures 6 through 11. The times mentioned in the description below are all tines after rupture. The abrupt, dcuble-ended rupture of a main steam line outside conta1==+2t causes a rapid loss of pressure in the steam generators. Figure 6 presents the outlet prescure of the affected steam generator. Since no credit was taken for backpressure in the steam piping downstream of the unaffected steam generator, the pressure transient in the unaff ected generator is almost identical to that shown in Figure 6 for the aff ected generator. A low steam w I jG a.i. I-12

pressure feedwater latching signal would be generated within 2 seconds in the steam headers. However, ro f eedwater latching was assu=ed. The safety grade main steam isolation valves would begin to close as a result of the latching signal, but do not close fact enough to have a significant effect within the ti=e frame of this analysis. Decrease in steam pressure lovers the saturation temperature and causes sene of the liquid in the steam generator to flash into steam, increasing the froth level in the stean generator. These eff ects cembine to increase the heat transf er from the primary coclant system. Main feedwater flew is asse=ed to increase to about 136% of rated flow at 2 seconds after rupture, and is assumed to re=ain constant thereaf ter. This high feeding rate tends to =aintain a high froth level in the steam generators. The core moderator average temperature decreases rapidly due to overcooling, as shown in Figure 7. The cooldown rate averages about SI/see over a period of 40 seconds. Core out m. coolant pressurt is shown in Figure 8. Pressure decreases very rapidly follewing rupture as primary coolant in the steam generators is overcooled. The pressure decrease is arrested =cmentarily when coolant in the highest elevations of the hot legs nears saturatica, making the system more compressible, and insurge frne the pressurizer begins to replace the primary coolant shrinkage. The low reactor coolant pressure ESFAS signal occurs at about 3 seconds, initiating high pressure injection at 8 seconds. System pressure decreases until the pressurizer beccmes e=pty at about 20 seconds. Voiding in the hot legs, upper region of the steam generators, and the reactor vessel upper head region act to slow the rate of pressure decrease. Core ma -' L ' > /G J' s I-13

9 9 flooding begins at about 25 seconds and also acts to slow the rate of system depressurization. The variable low pressure reactor ~ trip setpoint is trached at 0.4 seconds, and reactor trip occurs at 0.7 seconds. No turbine trip was assu=ed to occur. Total core reactivity is shown in Figure 9. After full control rod insertion at about 2.7 seconds, reactivity increases in the positive direction due to cooldown of the fuel and moderator. Boron injection frcc the core flooding tanks terminates the loss of subcritical margin at 28 seconds. Thereaf ter, boron injection f rcm the core flooding tanks, high pressure injection, and, if necessary, lew pressure inj ection is sufficient to increase the suberitical nargin as the fuel and moderator centinue to cool down. Figure 10 shows the response of total generated power in the core, which consists of fission power and fission product decay heat. The initial rapid decrease in system pressure following rupture causes a decrease in moderator density. This is due to the fact that a pressure reduction will reach the core a few seconds before the cooler fluid from the steam generators enters the core. This initial decrease in moderator density causes a decrease in core power. Core power decreases rapidly when control rod insertion begins at 0.7 seconds. After about 10 seconds, power level increases slowly due to loss of subcritical margin until core flooding boron injection terminates the loss of reactivity margin. Core power peaks at 28 seconds with a total generated power level of 14.1% f ull power, consisting of 9.3%F7 fission heat generation and 4.8% FP decay heat generation. w ~ I-14 ,i ,j e aa a

9 O The core inlet fluid enthalpy is shown as a function of ti=e in Figure 11. The total power, core outlet pressure, core inlet enthalpy, and core inlet flow rate calculated by TFA?2 were corrected f or control band errors and then fed into the RADAR ce=puter ccde for core ther=al-hydraulic analysis. 2. Reactor Coolant Flow Coastdown Case The input assc=ptions for the reactor coolant flow coastdown case were indentical to those for the full flow case, except tha t a four reactor coolant pc=p coastdown was asse=ed to begin coincident with rupture. A chronological sequence of events for this case is given in Table 4, along with the sequence of events for the full flev case. The response of key system variables is shown as a function of ti=e in Figures 12 through 18. The trends of the syste variables are the sa=e as in the full flow case, but the rates of pri=ary systes cooldoen and depressurization are slower due to coolant flow coastdown. A ce=parison to the full flow case will be =ade below. Two reactor trip signals are received at approximately the same time, on the reactor coolant pu=p =onitors and on variable low coolant pressure. The toolant flow coastdown reduces.the rate of heat transfer in the steam generators, which reduces the cooldown rate in c prt=ary syste=. Note in Figure 13 that the flow coast-down begins to significantly affect the cooldown rate at 15 to 20 seconds after rupture. Care flooding is not actuated until 52 seconds,'ter=inating the loss of suberitical reactivity =argin. After reactor trip, the generated core power level re=rins relatively constant below 10% full power. Power begins to increase slightly n 3 r-at about 55 seconds, but is ter=1nated at a very Icv level by yy baron injection from the core flooding anks. The cininum sub-critical reactivity =argin (See Figure 15) is s= aller than in the 1-15

full coolant flow case, pri=arily because the slower system cool-down allows the f uel to reach a lower temperature for any given moderator temperature. Since =oderator temperature is closely coupled to system pressure, a sicwer cooldown allows more time for fuel temperature decay prior to shutdown of the core by core flooding baron injectica. Even though the minieum subcritical margin is smaller, the additional tire for decay of r.he delayed neutron precursors causes the suberitical return to pcwer to be substantially less severe than in the full coolant flow case. Total generated core power peaks at 10 FP, consisting of 6: FP fission heat generation and 47. FP decay heat generation. The total power, core outlet pressure, core inlet enthalpy and core inlet flow rate calculated by TRAP 2 were corrected for control band errors and then fed into the RADAR computer code for core thermal-hydraulic analysis. w S j L: ..] o I-16

RESUI.TS OF THE RADAR ANALYSIS At the beginning of the stea= line break accident and prior to reactor scra=, the hot channel was assu=ed to have the =ax1=um design conditions as eu, tined above. At reacter trip, a fully withdratn control rod was assu=ed stuck out of the core and eff ecting this same hot channel. As the control rods were inserted, the channel telative pcwer density (RPD or FaH) was increased fro = 1. 8 to a value in the range of 6.0 to 9.0 and the axial flux shape in this channel was increased to a peak of 3.0 located a distance one third the length of the channel (Figure h,. The axial flux shape assumed after reactor trip (Figure li) is considered representative of the axial power distribution under local voiding conditions. ' i I-17 iU .) /v

O e = .E. =*.-- 6-b e e e

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Local voiding effecta on core reactivity and power distributions were discussed in Rafarence 1. Core reactivity effects associated with local voiding are nearly identical to thene associated with control rod perturbations.

Hence, for the situation of voiding in the upper half of the hot assembly, the axial power distribution is repres2nted by a,artially inserted control rod.

The axial peak of 3.0 results from a control rod insertion of 50% and represents the av.ial distribution for voiding in the upper half of the assembly. Two caces were examined for the return-to-power following a main steam line break accident. The first case sssumed the reactor coolant pumps would continue running thrcughout the accident transient, and the second case assumed that tne pumps tripped at the instant of steaa line rupture. The analysis consisted of running the accident several times on the RADAR computec code and varying the hot channel relative power density (RPD) af ter scram and during the return-to-power. This was necessary in order to achieve convergence with predictions of the -ari-um sustainable RPD for a given channel voiding. Figure 'O shows the =aximum sustainable channel RPD as a function of channel voiding. Figure 20 also shows the average hot channel voiding in the thermal-hydraulic =odel for various assu=ed char.nel RPD's with the pump running for the entire accident transient. Convergence of the ther=al-bydraulic and nuclear models was achieved for a RPD of 8.2 during the return-to-power. Figures 21, 22, and 23 depict the hot channel voiding, DNBR, and local heat flux as a function of channel length for au PPD of 9.0. A minimum DNBR of 2.090 (=odified Barnett correlation) resulted in the pu=ps running case for a channel RPD of 9.0 at the time of maxinun return-to-power. Figures 24 through 27 show the results of the case in which the pumps trip at s team line rup ture. Convergence of the Thermal-Hydraulic and Nuclear Models s. was achieved for a RPD of 6.5, and channel voiding, DNBR, and local heat flux as D3 ^ ~: 1-19 /U _'. ) v

a function of channel length are shown for a RFD of 7.0. A minimum D::0R of 1.927 (modified Barnett correlation) resulted in the pumps trip at rupture case for a channel RPD of 7.0 at the time of caxi=um return-to-power. Saturation conditions were reached in the hot channel for RPD', in the range of 5.0 to 6.0 for both cases analyzed. Voiding increased rapidly for small increases in peaking after saturation. For the case with pumps tripped at rupture, the average channel void went from 257. for a peak of 6.0 to 80% for a peak of S.O. This lends added credence to the fact that entre =ely high peaks cannot be sustained in the pressere and temperature ranges necessary to cause the return-to power phenomena. This entire analysis was perfor=ed using the maxi =u te=perature and pressure errors in the =ost conservative =anner, the maximus core flow maldistribution, =aximum design peaking including nuclear uncertainty, =axt=us manuf acturing errers on asse:bly axial flow areas, and hot channel factors en 1ccal pin heat flux and pin power rating. The system analysis was designed to maximize the return-to-power phencmena. Even with the above assumptiens, and using stuck rod assembly peaking clearly greater than the predicted - nim 50% margin is decanstrated to a minimum DNER of 1.3. Q l [ (I v.). 1-20

O IV. Su= mary and Conclusions A. Senry of Analyns Results A sum =ary of analysis results is presented in Table 5. These results indicate that a substantial margin on DN3 exists with or without reactor ecolant pump coastdown. Since the assumptions used in the accident analysis and core thermal-hydraulic analysis are based on worst case conditions, the results of an accident scenerio allowing some mitigation by non-safety grade equip =ent would be even less severe. B. Proposed Changes to Technical Specifications The TMI-2, Cycle 1 Tech Spec rod index limit curves have been revised to incorporate a 2% shutdown margin required for MSLB. The limits were calculated for the worst ti=e in life conditions (BOC) and are listed below: Power Minimum Rod Index (%) (WD) 102 143 50 132 15 117 0 6f The revised shutdown margin Tech S od index limits assure that at least 2 shutdown margin (K,ff .9 8) exists in the core at all times in the cycle for a reactor shutdown with the most reactive red stuck out of the core. Previous startup tests in B&W. reactors have confirmed the conservatiam of 3&W shutdown cargin calculations. The TMI-2 physics tests will provide mearured data to verify the adequacy of the B&W calcula-tio ns for TMI-2, Cycle 1. [% / (.' 'l V w I-21

C. Conclusion The analysis described above de=oustrates that the safety features existing on the TMI-2 plant ate adequate to sitigate the consequences of a suberitical return to power during a =ain steam line break. The fact that no depar',ure from nucleate boiling is predicted to occur insures that no fuel failure will occur as a result of suberitical return to power, even with a conservative power peaking factor applied to the stuck centrol rod fuel asse=bly. The plant Technical Specifications insure that a sufficient suberitical =argin can be =aintained for the steam line break accident throughout Cycle 1, and that suberitical return to power will be less severe than that de=enstrated to be acceptable in this analysir. ,e .L a.. I-22

9

References:

1. dand carried trans=1ttal on 9/8/77 to T. M. Novak, Chief, Reactor Systems Branch, NRC, providi g docu=entation on generic meeting of 8/16/77 regarding s tea = line b reak. mW ia q n.3 lU 's " e 1-23

Tab le 1 KEY ANALYSIS ASSUMPTIONS 1. Double-ended rupture of a main steam line outside centainment. 2. No credit for backpressure to unaffected steam generator from steam piping volume. 3. Bounding runout main feedwater flow rates. 4. 125% emergency feedwater initiated at 2 seconds. 5. No main feedwatur *. solation. 6. No main steam isolation for 2 minutes. 7. No turbine trip. 3. Initiated f rom rated full power; DNBR calculations based on 1C2% full power. 9. Fouled steam generator inventories. 10. Most negative end-of-cycle (Cycle 1) reactivity feedback parameters. 11. 2% ak/k shutdown margin at hot shutdown with most reactive control rod stuck out. 12. Offsite power available. 13. No credit to: primary system thick netal heat capacity. 14 Analyzed with mad without reactor coolant flow coastdown. 13. No hydraulic upsets of the condensate or main feedwater pu=ps. 16. 'a'orst cctdition of offsite power (available or unavailable). i t.l g I-24

TAELE 2. COMPUTER CODES USED TRAP 2 BAW-10128, AUt:UST, 1976 RADAR BAW-10069A, P'V.1, OCTOBER,1974 PDQ07 BAW-10117A, JANUARY, 197, e e"- g q I-25

6 TABLE 3. KEY IfFtX PA?#ETERS FOR iPAP2 POWER LEVEL, ?,Wt 2772 VARIABLE LOW RC PRESSURE FSAR FIG. 7.2-2 TRIP SETPOINT ECUIVA!lNT. DOPPLER FUEL -1.53x10-5 TE?.PERATURE COEFFICIENT, sK/K/F EQUIVALENT MODERATOR TE9ERATURE -3.0 x 10-4 COEFFICIENT, sK/K/F ASSU?.ED CONTROL R0D WORTH -4.23 AT 532F, I sK/K EOL SHUTDOWN P.ARGIN, % sK/K -2.0 INITIAL BOR0fl CONCENTRATION, PPM 17 NUMEER OF CORE FLOODIhb TAtlKS 2 CORE FLOODING ECRCN CONCiliTRATION, PPM REACTOR VESSEL TO CHECK VALVE 17 CHECK VALVE TO UPSTREAM CHECK VALVES 460 CORE FLOOD!tlG TANKS 2270 CORE FLOOCING TANK COVER GAS PRESSURE, PSIA 615 LOW RC PRESSURE ESFAS SETPOINT, PSIA 1615 NUMEER OF HPI PUPPS ASSUMED 2 ASSUMED HPI BORC:1 CONCEiTRATICN, PPM 17 eSSU?.E E5RGBlCY FEEDWATER FLOW RATE, GPM 1250 ASSUMED Mall FEEEWATE.' RUNCUT FLOW RATE, 15 TOTAL 3REAK \\REA, FT2 5.31 R)LLED STEAM E'E ATOR I?tEiTORY, L5i FER OTSG 4E, /U -.m I-26

TABLE 4. CHRONOLOGICAL SEQUENCE OF EVE'<TS T!hE AFTER RUPTURE (SEC) FULL RC RC FLOW EYE 11 FLOW C0ASTDOWN 1. RUPTURE OF 24" SCH, 60 0 0 MAIN STEAM LINE OUTSIDE CONTAINMENT 2. REACTOR TRIP ON 0.7 0.7 VARIABLE LOW PRESSURE 3. EMERGENCY FEEDWATER 2 2 INITIATED ~ 4. HIGH PRESSURE INJECTION 8 10 BEGINS (N0 BORON INJECTION) 5. CORE FLOODING BEGINS 25 52 6. MINIMUM SUBCRITICAL 28 65 MARGIN AND MAXIMUM RETURN TO POWER OCCUR =f R l v I-27

TAELE 5.

SUMMARY

OF RESULTS RC FLOW FULL RC C0ASiDOWN FLOW AT RUPTURE MINIMUM SUBCRITICAL -0.54 -0.32 MARGIrl, % 2K/K MAXIMUM RETURN TO 9.3 6.0 POWER (FISSION POWER), % OF 2772 MWt MAXIMUM RETURN TO 14.1 10.0 POWER (FISSION PLUS DECAY POWER), % OF 2772 MWt EOUNDING PADIAL X LOCAL 9 7 PEA'<ING FACIOR EOUNDING TOTAL PEAKIllG 27 21 FNTR MINIBUiDNER 2.09 1.93 I "3

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ATTACE EU 11 HOT CHANNEL THERMAL HYDRAULIC ANALYSIS FOR TdE STFAM LINE BREAK ACCIDENT I. INTRODUCTION The purpose of this report is to present the results of an analysis to determine the worst-case core ther=al hydraulic conditions during a main steam line break accident. A separate steam line break analysis has been sebmitted which dea's with the poteatial for fuel damage during a suberitical return to power with a concurrent abnor=al power dist ribution in the stuck control rod assembly. That analysis de=custrates that no departure from nucleate boiling (DNB) will occur as the result of a tuberitical return to power. However, in the performance of that analysis, we have identified more severe ther=al hydraulic results than those previously presented for the design hot channel analysis. Transient het channel DNBR results were presented in the FSAR, Appendix B, for the main steam line break with single active failures. The Appendix 153 resulto were calculiced using a combination B&W FLASEP and EDS RELAP4 compute ! model for tne blowdown analysis an'd the B&W RADAR computer crae for the core thermal hydraulic analysis. Appendix 13B concludes that no DNB will occur during the accident, with a =inimum DNBR of 1.7 for the worst case accident scenerio and worst-case core therma.1 Sydraulic conditio ns. Re cent calculations performed in conj unction with the analysis of subcritical return to power during the steam line break have indicated that the core thermal hydraulic results are inconsistent with those presented in Appendix 15B. Using the very conservative RADAR assumpt Lons employed for the Appendix 153 rhe umal hydraulic analysis, and using worst-case results from the recent TRA?2 accident an alys is, DNB is predicted to occur at reactor trip. II-l ql n1A /I k 8 I
The stess line break single f ailure analysis presented in the FSAR, Appendix 153, was designed to maximi:e mass and energy release rates to containment and the potential for cora return to criticality during cooldown. Feedwater flow rates were established using a coupled model. The FLASHP computer code was used to calculate the blowdown rate and the primary systen response. The RELAP4 computer code was used to calculate the main feedwater flow rate. In order to obtain agreement of results where the FLASHP and RELAP4 models overlapped in the steam generators, the steam generators were modeled in steady state with a disproportionate amount of inventory in the downcomer annuli, and an unrealistically low boiling height in the tube section. This approach was found to produce a more rapid decrease in steam pressure, which caused higher feedwater flow races. Thus, the coupled model, while unrealistic, was conservative for purposes of total mass and energy release and primary system over-cooling. In addition, a very large phase separation velocity was used in the steam generator tube section in the FLASHP model in order to prevent the release of two-phase fluid. This assumption is also conservative in terms of total energy release and total overcooling of the primary system. However, the combined assumptions of low tube region inventory and high phase separation velocity tend to depress the boiling height and give low primary to secondary heat trans fer rates in the firs t few seconds of the transient. Thus, the short term rates of pressure and temperature decrease in the primary system are no t as rapid as would occur if the iciling height increases rapidly following rupture. It has biten determined that n i,~ Q )I L,s II-2 /
hot channel DNBR at reactor trip is sensitive to the depressurization rate in the primary system. Tnereface, the hot channel DNBR results presented in Appendix 153 are not based on the worst-case core boundary conditions for the first few seconds of the transient. An analysis has been performed to eatablish the worst-case core conditions during the early stages of the steam line break accident. Kay assumptions and resulta for this analysis will be presented la subsequent sections. The 3&W TRAP 2 computer code was used to calculate the blowdown and primary system response. It should be noted that the non-conservative nature of the DN3R results presented in Appendix 13B was not due to use of the FLASHP code per se, but rather due to the failure of user models to provide worst-case core conditions in the initial stage of the blowdown. The multi-region steam generator heat transfer mocel available in the TRAP 2 code allows core realistic simulacion of transient heat transfer. This reduces the range of uncertainty which would have to be bounded by conservative s1=plifying assumptions if the FLASHP code were applied. The results of the TRAP 2 calculation were used to establish the core boundary conditions input to the S&W RADAR computer code. RADAR was used to calculate transient DNBR and fuel thermal conditions. II. DESCRIPTION OF BOUNDING CASE A conservative scenario has been established which =aximizes the adverse effects of the early portion of a steam line break on hot channel DNBR. This case is very similar to those presented for the analysis of core conditions during the subcritical return to power. = c.u 77,3
The accident is initiated from full power, which is the most adverse pcwer level with respect to DNB, and also provides the greatest steam generator inventory available for overcooling. The as s ume d double-ended rupture of a main s team line, maximum runout feedwater flow rates, and early initiation of emergency feedwater all act to maximize the rate of heat transfer f rom the primary sys tem. This results in the maximum rate of depressurization of the primary coolant. ? The assumed total break area of 7.66 f t~ co rresp onds to the double-ended rupture of a 29 1/4-inch s chedule 80 non-seismic Category I line. However, 24-inch sections upstream and downstream of the break location limit flow to the break. The 8-inC turbine bypass cross-piping provides an additional flow path to the break. The total ef fective area for choked flow to the break location is only 5.633 f t In the TRAP 2 calculations performed for the core return to power analysis, the initial rate of primary coolant depressurization was so rapid that the resulting decrease in core =aderator density provided negative reactivity feedback at the end of cycle. This caused a reduction in core power level prior to reactor trip. In order to maximize the power level prior to trip, beginning of cycle reac-tivity feedback parameters were used in this analysis. In order to maximize the reduction of reactor coolant flow, a four RC pump coast-down was assumed to begin coincident with rupture. A su==ary of cey assumption is presented in Table 1. Ill. RESL'LTS OF mE ANALYSIS A. TRAP 2 Results The transient responses of key parameters which af fect DNBR are shown in Figures 1 th rough *. Note that r"oture occurs at II-4 gl [i i ~/ it u s
~~I second in these figures.~~
Following rupture, che core coolant pressure and enthalpy decrease rapidly. Core inler flow rate decreases due to a four reactor coolant pump trip at rupture. The average core moderator density decreases initially due to depres s urization. At beginning of cycle, this causes a slight increase in fission rate (see Figure 1). The reactor trips on the reactor coolant pu=p =enitors with a 0.65 second trip delay. The power generation rate decreases rapidly as the control rods are inserted. Coolant pressure decreases rapidly until the fluid in the highest elevations of the hot legs nears saturation conditions, then pressure levels out momentarily, with surge flow f rom the pressurizer corpensating partially for shrinkage of the primary coolant. About 5 seconds af ter rupture, coolant pressure begins to decrease again. The para =eters plotted in Figures 1 through 4 were fed into the RADAR code for core ther=al-hydraulic analysis. ,l 0u.v II-5
B. IV.DAR Recults The thermal-hydraulic evaluation of the steam line break accident transient used the closed channel transient RADAR code (SAJ-10069A, Rev. 1). Two sub-channels were modeled into the RADAR code. The first sub-channel represents an average core channel and it used to calculate the core pressure drop as a function of time during the transient. The core pressure drop, as calculated by RADAR, agrees wi th the s teady-state codes in which the en ti re core is modeled. The input to RADAR consists of system flow, core pcwer, system pressure, and ccolant inlet temperature as a function of time. The core pressure drop versus time was applied across a second channel to calculate the r.?cimum DNBR, clad temperature, fuel tempera ture, etc. This second channel assumed the wors t nuclear, thermal and mechanical conditions exist simultaneously. If this particular channel meets all the thermal design criteria, then it can be safety assumed that all other channels in the core will be no worse thermally than this limiting channel. The maximum design conditions are represented by the follcwing assumptions. a. The !!mi ting fuel assembly (radial peak of 1.69 as shown in Figure 5) possesses the fuel pin wi th the maximum value of F (1.78) as determined 6g from examination of the reference desig, radial peaking distribution and frem examination of the maximum, ncminal, ar.d minimuu fuel assembly spacing to find the effects en the local peaking distribution. i v,, II-6
b. The maximum value of F nuclear (max / avg axial fuel rod heat ihrut) is determined for the limiting steady-state conditica such that su f fi cien t ma rg in is provided to accommodate asymmetrical axial power shape considerations which will permit core operation with both poritive and negative power imbalance. The maximum errors on core pressure and inlet temperature are applied in the most conservative manner. They are -65 psia and +2F, respectively. Every channel in the core is assumed to have the nominal pressure drop c. associated with the core flow conditlans. d. The limiting fuel assembly is penalized 5% in isothermal flow such that it receives only 95% of the flow associated with an average core bunuie at 100% of system flow (four pump operation). The limiting fuel assembly is assumed to have a reduced peripheral flow e. area due to the proximity of the adjacent fuel assemblies. f. On the subchannel types wi th m.aximum values of F I** aH ' channel factors are applied : 1. The channel is assumed te have a reduced flow area, represented by the het channel factor F,, less than 1.0. 2. The fuel pin will have the greatest beat output oy vi rtue of the F ", 9 which increaser the local surface heat flux. 3 And F, which increases the overall fuel pin power rating, will be greater than 1.0. Severa. accident transients were investigated with the most limiting case resulting at beginning-of-cycle (BOC) and the RC pumps tripped at the instant of steam line rupture. The results of this analysis are shown in Figure 6 as transient responses of clad temperature, fuel temcerature, and minimum DNBR. ~ Although the min imum DNOR is shown to recover racidly, nc recovery from film boiling was assumed for the clad and f uel temperature calculations. Ql f'. _ i ' c - v II-7
No fuel damage will er : ult during this accident transient becausa there is always sufficient margin between the actual pressure di fferential across the cladding (SP) and the AP required for structual failure of the clad. Figure 7 shcws the relationship of cladding stress and failure temperature currently used in the ECCS evaluation model (BAW Tepleal 10104). Figure 8 shcws the same curves af ter converting the clad stress to SP from Ir. side to outside of the cladding. The most critical combination of pin pressure, clad temperature, and system pressure occurs in the range of 4 to 8 secs. af ter steam line rupture. From Figure 8, using the envelope of all data points and the maximum clad temperature, a SP across the clad of approximately 350 pst is required for failure. Figure 4 shows the system pressure during the transient. In the range of 4 to 8 secs, system pressure is between 1678 psia and 1603 psia. The highest pin pressure o curring anytime during cycle 1 operation is 1600 psia. Thus, there is 350 to 400 ps! margin to clad failure during the most critical time of the transient. For times after 8 sees, pin pressure and clad temperature are declining at such a rate that margin increases r.pidly even though there is a steady slew decline In system prsssure. These calculattens are extremely conservative and are intended to shew the margin available with the worst pessible assumptiens a. d do not reflect the true margin to actual condittens. The pin pressure used for these calculattens (1600 psi) would cccur only during the inftlel cycle start up before any fuel densificatien occurs and would never be exceeded during the first cycle. The RADAR cede model assumed maximum fuel densification to maximize the stored heat In the fuel ecds to obtain the worst case clad and fuel t empe ra tu re s. When the fuel is fully densified, pin pressure will be at a minimum. No credit was taken for the fact that maximum pin pressure and maximum fuel temperature are mutually exclusive events. Analyses Indicate thet very little, if any, of the ccre would have a CN8 If most probable scre assumptICns were used instead of tha maximum Dl ,o " 1 /~ b L i II-8
design parameters that were cutlined earlier. Figure 9 shows the MDNBR obtained during the transient as a function pin peaking and indicates that, at =ost, 15% of the core would experience DNB. These calculations again are very conservative and assume maximum temperature and pressure errors, design peaking distribution, and initial 102: core overpower conditions. The effect of abnormal peaking conditions in a stuck control rod fuel assembly was investigated in relation to DNB at reactor trip. As mentioned earlier, DNBR in the design hot channel recovers very rapidly following reactor trip, even though no credit was taken for recovery from DNB in the clad and f uel temperature calculations. If a stuck rod peaking distribution is i= posed on the design hot channel following reactor trip, the elevation of minimus DNBR shif ts to a location one-third the length of the channel. As demonstrated in a separate report on suberitical return to pcwer during steam line break, no DNB occurs as a result of return to power even in the stuck control rod fuel asse=bly. In addition, the shif t in power distribution to the bottom of the fuel assembly causes peak clad temperature to be lower than that shown in Figure 6 for a fuel pin which enters DNB at reactor trip. Thus, subcritical return to power with abnormal peaking in the stuck control red fuel assembly will not aggravate the clad and fuel temperatures presented for the above analyses, and DNE at reactor trip can be analyzed as an independent phenomenon. IV. Conclusions The very conservative analysis presented above demonstrates that fuel f ailure will not occur as a result of adverse core conditions in the initial stage of a steam line break during Cycle 1. Even for wo rst-case design assu=ptions, there is a significant margin on cladding ap required for clad failure 9 I (i "' 9 II-9 ~, ' L t_ c
and on clad temperatures which would create a concern with metal-water reaction. A separate report has demonstrated that no DNB, and therefore, no fuel failure, occurs following reactor trip. We cotclude that core performance during the steam line break accident will be acceptable throughout Cycle 1. The results of this analysis supercede the results of the D iBR analysis presented in the FSAR, Appendix 15-B. No other results or conclusions presented in Appendix 15-B are ac= ended by this report. Gl l' ' ' ? 11-10 /i s
Table 1 Initial Conditions and Assumptions Rated Power Level, MWt 2772 MWt Initial Power Level for TRAP 2 Analysis, 100 % Rated Initial Power Level for RADAR Analysis, 102 % Rated -5 Equivalent Doppler Fuel Te=perature -1.25x10 Coefficient, ak/k/F Equivalent Moderator Temperature +0.5x10 ' Coefficient, ak/k/F As s ume t. Control Rod Worth at -2.9 532F, % ak/k BOL Shutdown Margin, % ak/k -2.0 Pump Monitor Reactor Trip Delay Time, sec. 0.65 Assumed Emergency Feedwater Flow Rate, gpm 1250 Assumed Main Feedwater Runout Flow Race, % rated 136 Fouled Steam Generator Inventory, 43000 lbm per OTSG Total Break Area, ft 7.66 Effective Choked Flow A;ea for Available Flow 5.633 2 to Break, Ft M II-11 mi n ',1 ( [. 8
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ML19220C332

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